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StRUCtURAL MonItoRInG And SEISMIC AnALySIS of A bASE-ISoLAtEd bRIdGE In doGnAC. Bedon1, A. Morassi2

1 Department of Engineering and Architecture, University of Trieste, Italy2 Department of Civil Engineering and Architecture, University of Udine, Italy

Introduction. In recent years base isolation has become an increasingly applied structural design technique to protect bridges and buildings from severe earthquakes. Main goal of base isolation is to produce a substantial decoupling of the superstructure from the substructure resting on the shaking ground, minimizing the internal state of stress during an earthquake by increasing the period of the structure and by acting as energy dissipation device (Naiem and Kelly, 1999). This structural solution also has the benefit of giving better distribution of the seismic forces between the various elevation supports. In the specific field of bridges, base isolation techniques were initially confined to long structures, but in several countries – due to more severe seismic codes – base isolation has been recently applied also to small and medium-size bridges. In this paper, results of an experimental/analytical study focused on the dynamic behavior of a 75m long, reinforced concrete (RC), post-tensioned, based isolated bridge are discussed.

The bridge was built in the Municipality of Dogna (Friuli Venezia Giulia), in an area with high seismic activity. The structure replaces an existing bridge – positioned a hundred of meters upstream – that highlighted marked hydraulic inadequacy during an exceptional flood (August 2003). Within the works carried out to restore the Fella river hydraulic regime, the new bridge reduced occupation of the river bed (e.g. single pier support). The mountain environment, moreover, suggested the use of a slender and low-impact structural solution, although with appropriate seismic resistance, hence preferring the base isolation technique (Alessandrini et al., 2009) to traditional supports (e.g. fixed or unidirectional bearings). In May 2007 the bridge underwent an extensive series of static truck-load and harmonic forced vibration tests with low levels of excitation. One of the purposes of the investigation was to verify the reliability of finite element (FE) models to describe the measured dynamic behavior of Dogna bridge. Dynamic data (namely, natural frequencies, mode shapes, and damping factors) can in fact provide meaningful results, if used to improve the FE model of a bridge, enabling for example to properly estimate mechanical properties of materials and boundary conditions. Another important purpose of the present research was to define a baseline model of the bridge for future diagnostic investigation, being this issue is of great importance for the company who manages the local highway network. A long-term plan of maintenance of the bridge, in this sense, could not neglect possible changes of performances of the bearing devices (Bonessio et al., 2012). Even though their response variations have been the goal of research at material and device level, procedures for monitoring the device functionality in service should be in fact properly developed. The periodic removal of isolators from bridge for a laboratory test campaign, for example, would not appear as a feasible approach, due to high economical impact. A more reliable solution, otherwise, could be represented by periodical testing of the bridge with devices in service, to localize and quantify global response changes generated by local degradation of structural components, as well as by isolation devices. Most of the diagnostic methods for damage detection in bridges are in fact based on comparison between the structural responses of the actual - possibly damaged - state and the reference (undamaged) configuration of the bridge. It follows that, to give aproper interpretation of the damage-induced changes on its dynamic characteristics, it is crucialto have at disposal an accurate knowledge of the undamaged configuration of the bridge, andthis was another important goal of the present study. In this paper, experimental and analyticalinvestigations are recalled and summarized from a recent extended research study (Bedon andMorassi, 2014). Harmonically forced tests performed on Dogna bridge are used for extractionof its dynamic properties, via experimental modal analysis (EMA) techniques. An accurate 3D

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FE-model is developed and FE model updating is carried out to reduce discrepancies between analytical and experimental dynamic estimations. The seismic behavior of the bridge is then predicted using suitable ground motion records and an optimized, computationally efficient, 2D FE-model that could represent the baseline model for future monitoring and condition assessment programs on Dogna bridge.

The Dogna Bridge. Dogna Bridge is a two span, two-lane post-tensioned reinforced concrete (RC) bridge.

Construction of the bridge, designed in accordance with NTC2008, was completed in Spring 2007. Main features and nominal dimensions are shown in Fig. 1a. Each span is 37.5 m long and a single RC pier (2.4 m thick, 4.0 m deep and 10 m high) is present at mid-span. Pier was built on cast-in-place concrete, 18 m long piles (diameter 1.2 m). The lateral abutments, consisting of vertical RC walls, are supported on 22 m long cast-in-place concrete micro-piles (diameter 0.3 m). At mid-span, the deck is continuous and simply supported – both on the abutments and on the pier – by two multi-directional elastomeric bearings (Fig. 1b). The bearings are seismic isolators (type SI-N-1200/112, FIP Industriale) with 1200 mm diameter, made up of layers of steel laminates and hot-vulcanized rubber. For large shear strain, their compound is characterized by nominal dynamic shear modulus Gdin= 0.8 MPa and equivalent viscous damping coefficient ξ= 10-15%. The bridge cross section is uniform along the longitudinal direction, except near the pier. The typical cross-section consists of a 5.0 m × 1.3 m central rectangle, which contains the prestressing tendons, and two lateral portions with thickness varying from 0.5 m to 0.2 m. Near the pier, the thickness of the deck cross-section gradually increases from 1.3 m to 2.2 m (Fig. 1a).

Dynamic testing. In May 2007 the bridge underwent an extensive series of harmonic forced vibration tests with low levels of excitation for determining the frequency response function (FRF). Steady-state harmonic vibrations were induced by means of the stepped-sine technique. A unidirectional mechanical actuator with out-of-balance rotating masses furnished a sinusoidal load acting along a given direction and with amplitude proportional to the square rotating

Fig. 1 – Dogna bridge: a) plan, elevation and transversal cross-sections (on the pier and at the centre of a single span); b) seismic isolators: cross-section, overview and nominal characteristic behavior.

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frequency. The mechanical exciter, capable of generating forces with maximum amplitude up to 20 kN, was mounted on the bridge deck at one third of the span on Dogna side, close to the downstream sidewalk (Fig. 2a). The acceleration response of the bridge was measured by seismic accelerometers (ICP Series 393B12) with sensitivity equal to 10 V/g. During the tests, the rotating frequency of the actuator was gradually increased in the ranges 1-6 Hz and 6-15 Hz. A data acquisition unit was used for data acquisition, control of the force actuator and automated signal processing. Structural response was simultaneously measured by 18 sensors (16 placed on the bridge deck and 2 on the top section). 8 accelerometers with vertical axis were mounted on both the sides of the deck, e.g. at the quarters of the right span (side Dogna), and at mid-span of both the spans. Longitudinal and transverse motions were monitored at the mid-span cross-section of both spans (Fig. 2a). Based on the experimental layout proposed in Fig. 2a, the response of the structure was separately measured under excitation parallel to the vertical (V), longitudinal (L) and transverse (T) direction. Tests were carried out in absence of the deck asphalt layer. FRF measurements were carried out with a frequency increment of 0.05 Hz. A smaller step (0.02 Hz) was used in the bands 1.8-3.4 Hz and 6.5-8.5 Hz, to improve the identification of vibration modes under vertical excitation.

Fig. 2 – Dynamic experiments: a) instrumental layout; b) example of comparison between measured and synthesized point inertance; c) identification of EMA Modes 6 and 7; d) 3D view of experimental modal shapes.

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Interpretation of dynamic test results. Careful analysis of experimental FRF curves suggested the use of responses to excitation along the transverse (T), longitudinal (L) or vertical (V) directions for the reconstruction of the dynamic properties and vibrating modes with prevailing amplitudes along the T, L or V-directions respectively. The experimental FRF curves resulted almost regular in the low frequency range (1-6 Hz). Some difficulties have been encountered in the analysis of FRFs at high frequencies (12-15 Hz), due to the presence of more pronounced irregularities and low amplitude measurements. Fig. 2b shows a comparison between some measured and synthesized point inheritance functions of the bridge.

Vibration modes generally resulted in resonance frequencies well separated, with the exception of the pairs of modes (2, 3) and (6, 7). Modes 2 and 3 were easily identified, since their modal shape corresponds to vibrations with prevailing amplitudes in the horizontal plane and along the vertical direction, respectively. A different situation was encountered for Modes 6-7, since both of them are primarily vertical modes. In this case, it is found convenient to analyze the FRFs obtained as the half-difference and the half-sum of the FRFs measured in points belonging to the same cross-section of the deck. In general terms, the half-sum allows in fact to remove torsional contributions in the measured deformations, thus to take into account the bending vibrations only. Conversely, the FRF obtained as half-difference emphasizes the torsional component of the modal deformation (Fig. 2c). This method allowed to reconstruct separately Modes 6 and 7, under the assumption that amplitudes of oscillation of the points belonging to the transverse cross-section of the deck are equal (in absolute value). Tab. 1 summarizes the results of the Experimental Modal Analysis (EMA). Natural frequencies are the average values obtained from the analysis of FRF curves. As shown, deviations from the average values can be considered negligible for modes at low frequency, while differences generally increase with the mode order. The estimation of the damping ratios is equally good, although differences are important for some higher order modes. Based on damping ratios collected in Tab. 1, it can be observed that bending and torsional modes of the deck generally have lower damping terms (≈ 0.7-1.8%), compared to typically high damping values of almost-rigid body motions (≈ 2.9-4.3%), in which the energy dissipation arising from the elastomeric bearings can be appreciated.

Tab. 1 - Experimental Modal Analysis results. Mean value of natural frequency f and damping ratio ξ, with corresponding maximum deviations. B= Bending; T= Torsional; RB= almost rigid-body motions.

r [-] Description [-] f [Hz] ξ [%]

1 1st B 2.022 ± 0.001 0.88 ± 0.03

2 RB Transverse 3.053 ± 0.003 2.88 ± 0.05

3 2nd B 3.180 ± 0.002 0.89 ± 0.05

4 RB Longitudinal 3.�05 ± 0.002 4.33 ± 0.07

5 RB Torsional 4.831 ± 0.011 3.93 ± 0.13

� 1st T �.887 ± 0.04� Not available

7 3rd B �.924 ± 0.015 1.15 ± 0.20

8 2nd T 7.995 ± 0.005 0.88 ± 0.10

9 4th B 9.107 ± 0.020 1.78 ± 0.44

10 Coupled B-T 12.910 ± 0.025 1.�� ± 0.20

11 Coupled B-T 14.228 ± 0.081 0.�� ± 0.12

12 Coupled B-T 14.433 ± 0.100 0.77 ± 0.27

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The presence of the elastomeric supports on abutments and pier typically resulted in decoupling of the dynamic response of the bridge. EMA modes collected in Fig. 1d can be in fact grouped into two classes, namely vibration modes corresponding to (almost) rigid-body deformation of the deck and vibration modes associated with vertical flexural/torsional deformation of the deck. The first class includes EMA Modes 2, 4 and 5, which correspond to a rigid transverse (3.05 Hz), longitudinal (3.61 Hz) and torsional (4.83 Hz) modes, respectively. Concerning the other modes, Mode 1 has a natural frequency of 2.02 Hz and corresponds to the fundamental bending mode of a two-span beam. In it, torsional effects of the deck are almost null and the vertical modal components evaluated on a pair of transversally aligned points differ from each other by about 10%. The two spans show in-phase vibrations in Mode 3 (3.18 Hz), with spatial shape comparable to the second bending mode of a simply supported continuous two-span beam. A loss of symmetry in the longitudinal direction emerged in this mode, with amplitude of vibration at mid-span of Dogna side which is about 20% larger than the corresponding amplitude

Fig. 3 – Numerical simulations: a) overview of 3D FE-model; b) reference model for the solution of the eigenvalue problem (3D-REF); c) 2D FE-model, cross-section on the pier; d) 2D FE-model, correlation with static truck-load experiments (flexural load scheme; longitudinal cross-section of the deck); e) example of a group of accelerometers.

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measured on the other span. Mode 7 (6.93 Hz), is reminiscent of the third bending mode of a two-span continuous beam, although the limited number of measurement points on the span opposite to Dogna side made difficult the complete description of the modal shape. Mode 6 (6.89 Hz) and Mode 8 (8.00 Hz) are representative of the fundamental and second torsional modes of the bridge deck. A significant coupling between bending and torsional deformation of the deck was finally detected for the higher four EMA Modes 9-12.

FE-model calibration. A 3D FE-model was developed with the SAP2000 computer package, based on the nominal geometrical details of the bridge (Fig. 3a). The typical FE-model was carried out using 3D solid FEs to model the deck and the pier. The inertial/stiffening effects of footways were neglected, while the abutments were considered as non-deformable (e.g. rigid foundations) and the pier was assumed to be rigidly connected to ground, hence neglecting possible soil-structure interaction effects on the dynamic behavior of the bridge. Concrete was described as a linear elastic, isotropic material. Based on experiments performed on 150 mm diameter cylinders casted during construction of deck slab and piers, Young’s modulus was set equal to Ec= 43.2 GPa, being νc= 0.15 and ρc= 2300 kg/m3 the Poisson’s ratio and density. Each seismic isolator was also described in the form of 3D solid elements, by taking into account the mechanical properties of an equivalent, orthotropic, linear elastic material able to provide – in the form of well-calibrated Young’s and shear moduli EX, EY, EZ and GZ, GY, GZ – the desired shear KX, KY and axial KZ stiffnesses for the devices. Different 3D FE-models partly discussed below (e.g. 3D-NOM, 3D-REF, 3D-OPT) were in fact characterized by the horizontal elastic stiffness of seismic isolators, being the geometrical and mechanical description of the other bridge components identical for all of them.

Preliminary 3D FE-model (3D-NOM). In the first case, preliminary FE-modal simulations were developed by taking into account the nominal shear stiffnesses KX and KY given by the producer of seismic isolators (KX= KY= 89.7 MN/m, based on Fig. 1b), with KZ= 7631 MN/m the corresponding nominal stiffness in the vertical direction. Modal analysis results are collected in Tab. 2. As shown, discrepancies up to 28% were found between analytical and experimental predictions, for the first ten vibration modes. Further studies and refined FE-models were consequently taken into account.

Refined 3D FE-model (3D-REF). For the 3D-REF model, a rigorous analytical method was taken into account for a proper calibration of seismic devices. The stiffnesses KX and KY, specifically, were separately estimated based on some experimental frequencies collected in Tab. 1 and simple analytical models able to properly describe the expected behavior of the bridge deck. In the longitudinal (Y) direction, in particular, the deck was rationally assumed to behave as a rigid-body simply supported on the abutments (with 2Ky the stiffness provided by two isolators, on each abutment) and the mid-span pier (with Ky

C the stiffness contribution given by two seismic devices working in series with the concrete pier), being:

and . (1), (2)

In Eq.(2), H is the pier height; A and χ are representative of the area and the shear factor of the pier cross-section; Iy its moment of inertia respect to the bending axis and Gc the shear elastic modulus of concrete. Assuming for the longitudinal mode of the bridge (RB Longitudinal, Tab. 1) a rigid-body motion, the sum Ky,TOT of stiffness contributions offered by the abutments and the pier supports should in fact result in:

, (3)

with Mdeck the total mass of the deck plate, hence leading, based on Eqs.(1) and (2), to:

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. (4)

Concerning the transversal (X-direction) stiffness of the same devices, due to transverse deformability of the deck in the horizontal plane suggested by the measured deformed shape of EMA Mode 2 (Fig. 1d), a different analytical approach was taken into account. The bridge deck, specifically, was described in the form of a two-span continuous beam having a uniform cross-section along the total deck length and resting on three elastic supports (Fig. 3b). In this hypothesis, the elastic stiffness of the end supports is in fact reasonably given by two seismic isolators only (Kx

a= 2Kx), while the transversal stiffness offered by the mid-span “pier-isolators” support should be calculated by taking into account both the contribution of two isolators and the deformability of the pier along the transverse direction:

. (5)

While the pier stiffness contribution pierxK was calculated in accordance with Eq.(2), the

transverse in-plane free vibration response of the deck (with fundamental frequency ω and modal amplitude u= u(y)) was estimated by means of a well-defined eigenvalue problem. In doing so, the moment of inertia I of the deck was assumed constant and coincident with the nominal value of its mid-span cross-section (Fig. 1a). Solution of the eigenvalue problem in closed form allowed then to estimate the stiffness parameter KX as the optimal value able to provide full agreement between the fundamental analytical frequency ω and the corresponding experimental frequency EMAf2 (Tab. 1).

Application of the aforementioned methods to the studied bridge resulted in KY= 172.4 MN/m and KX= 151.2 MN/m for the longitudinal and transversal directions respectively. An average value KX= KY= 161.8 MN/m was then reasonably assumed in the 3D-REF model. The isolator stiffness in the vertical direction, finally, was kept equal to its nominal value (KZ= 7631 MN/m). Compared to the 3D-NOM model, refined calculation of the shear stiffnesses KX and KY resulted in marked improvement of correlation between analytical modal predictions and experimental data (Tab. 2), hence justifying the discussed analytical assumptions and suggesting their possible applications to similar bridges. Maximum discrepancies between analytical and test frequencies generally resulted lower than 7%.

Optimal 3D FE-model (3D-OPT). Further improvement of the 3D-REF model was finally obtained by sensitivity numerical studies performed by means of parametric modal investigations. The shear stiffness KX = KY of each isolator, in this sense, was progressively increased – starting from the aforementioned 3D-REF value – so that discrepancy between EMA frequencies and corresponding 3D-OPT predictions could be minimized. The optimized value KX

OPT= KYOPT=

194.2 MN/m was identified as the stiffness value able to provide the best correlation between analytical predictions and test measurements. The obtained discrepancies were in fact generally lower than 3% (Tab. 2). In this case, it is interesting to notice that the optimal KX = KY was detected to be ≈ 2.2 times the nominal value suggested by producers (Fig. 1b), hence confirming the importance of a proper calibration of FE-model components for structural investigations or monitoring programs.

The equivalent 2D FE-model of the bridge. Based on accurate model updating of 3D models, a 2D-version of the 3D-OPT model was successively developed in SAP2000, to increase its computational efficiency but preserving its original accuracy (Tab. 2). The deck was described by means of shell elements belonging to two different horizontal planes, located at the barycentre of the central part of the cross section and of the two lateral portions respectively (Fig. 3c). Discrete Kirchhoff, isotropic, four node shell elements, with 6 DOFs at each node were used. The kinematical continuity conditions between shell element nodes

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belonging to different planes were then established by means of internal rigid-body constraints. The average dimensions of shell elements was comprised between 0.25 m - 0.50 m in the mid part of central span, whereas a refined mesh pattern was used in the regions close to the pier and to the lateral supports. The thickness of these shell elements was defined based on the actual longitudinal profile of the deck (Fig. 1a). The same shell element type was adopted also for the pier. The mass contribution offered by footways was then taken into account in the form of lumped masses. Concrete was mechanically characterized as for the 3D models. Each isolator, finally, was described in the form of three linear elastic springs applied at the intersection point between the vertical axis of the real isolator and the extrados of the deck. These springs, reacting separately along vertical, transverse and longitudinal directions with stiffnesses KZ, KX and KY respectively, were calibrated in accordance with the 3D-OPT model. In addition, two elastic rotational springs reacting in the longitudinal plane y-z and in the transverse plane x-z were placed at the constrained points, to take into account the elastic contrast offered by isolators. The stiffness Kφ of these rotational springs was evaluated as Kφ= πR4/4 KZ, with R= 0.60 m signifying the nominal radius of the isolator basis (Fig. 1b). Dynamic simulations highlighted that the calibrated 2D FE-model (2D-OPT) of the bridge reproduces well the vibration modes of the 3D-OPT Model. Discrepancies on natural frequencies generally resulted negligible for quasi-rigid body motions and for bending vibrational modes, with errors lower than 2%. Larger deviations (≈4-6% ) were found for torsional modes. In any case, the 2D-OPT model was considered sufficiently refined to perform further investigations on the bridge. Accuracy of the 2D-OPT model was suggested not only by agreement with experimental frequencies, but also by additional validation performed against static truck-load tests (e.g. Fig. 3d).

Seismic analysis of the bridge. In order to appreciate the benefits of the isolated system under earthquake motions, the seismic response of the isolated bridge was finally analyzed (2D-OPT) and compared to the dynamic behavior of the same structural system not equipped by seismic isolators (2D-FIX). In doing so, a bilinear constitutive shear force-shear displacement along the X and Y directions was used for the mechanical characterization of seismic devices (Fig. 1b), by

Tab. 2 - Correspondence between experimental (EMA) and numerical mode shapes of the preliminary (3D-NOM), refined (3D-REF) and optimal (3D-OPT) FE-models. Natural frequencies f and errors. ∆= 100×(= 100×(fr

(3D) – fr (EMA))/fr

(EMA). r= mode order; w.c.= without correspondence.

EMA 3D-NOM 3D-REF 3D-OPT

r f r f Δ MAC r f Δ MAC r f Δ MAC [-] [Hz] [-] [Hz] [%] [%] [-] [Hz] [%] [%] [-] [Hz] [%] [%]

1 2.022 1 2.045 1.12 99.4 1 1.985 -1.84 99.4 1 1.993 -1.4� 99.4

2 3.053 2 2.442 -20.01 93.� 2 2.857 -�.41 87.3 2 2.989 -2.10 85.3

3 3.180 4 3.292 3.53 89.� 3 3.1�5 -0.48 89.5 3 3.171 -0.29 89.�

4 3.�05 3 2.�82 -25.59 100.0 4 3.3�9 -�.5� 100.0 4 3.�09 0.10 100.0

5 4.831 5 3.�9� -23.49 95.9 5 4.�11 -4.5� 94.3 5 4.918 1.81 93.4

� �.887 7 7.�5� 11.17 85.1 � �.�05 -4.09 83.� � �.�84 -2.95 83.8

7 �.924 8 7.105 2.4� 98.0 7 �.809 -1.80 97.5 7 �.824 -1.59 97.5

8 7.995 10 10.254 28.2� 89.7 9 8.829 10.43 89.0 9 8.833 10.48 89.0

9 9.107 9 9.3�1 2.79 83.5 10 8.903 -2.24 73.7 10 8.90� -2.20 73.8

10 12.910 w.c. - - - w.c. - - - - - - -

11 14.228 1� 19.581 37.�2 53.3 15 1�.488 15.88 5�.5 15 1�.507 1�.02 5�.5

12 14.433 w.c. - - - w.c. - - - - - - -

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taking into account the identified optimal value of elastic stiffness (KX= KY= 194.2 MN/m) and the inelastic mechanical behavior provided by the manufacturer (Fig. 1b). An elastic link with stiffness Kz = 7631 MN/m was used along the vertical Z-direction. The structural solution of the bridge with traditional constraint scheme, in contrary, included two supports on each abutment and two supports on the pier. All the supports were located exactly in the same position of the isolators. Seismic analyses were developed according to the Nonlinear Time-History Analysis implemented in the software package SAP2000 and suggested by NTC2008. Seven groups of earthquakes – inclusive of X, Y, Z components – were taken into account. Earthquakes were defined by synthesized accelerograms randomly selected and representative of ground motions expected in the area of the bridge. Fig. 3e shows an example of the acceleration time-histories and corresponding pseudo-acceleration response spectra for the usual 5% structural damping, soil type B (average velocity of shear waves Vs,30 comprised between 300 and 800 m/s), peak ground acceleration ag= 0.35 g and importance factor 1.3 (e.g. construction of strategic relevance). Experimentally identified damping ratios were considered in the structural analysis (Tab. 1). A time length of 50 s and a step-resolution of 5×10−4 s were used in the numerical analysis. Main results are summarized in Tab. 3, in the form of maximum values of support reactions at the abutments and at the pier (average value of maximum reactions obtained for each group of ground motions). Reaction forces are proposed for the optimal 2D-OPT model and the traditionally restrained 2D-FIX models. Based on collected predictions, it is clear that the global seismic force transmitted by the superstructure to the elevation support components is about the 30% of the corresponding value for the bridge with traditional constraint setting. In addition, base isolation system allows to distribute seismic force effects an all the vertical resisting elements, avoiding the concentration of forces on a single structural component. Maximum shear displacements of isolators under earthquake resulted equal to 72 mm and 84 mm along the X and Y directions respectively. Finally, as expected, dynamic analysis showed that the incursions of the isolators into the inelastic range during ground motions are significant, and that these devices are responsible for the energy dissipation through the hysteretic behavior and for rising the period of the lower vibrating modes of the bridge. In Tab. 3, it can be also seen that improper mechanical calibration of seismic isolators would provide inaccurate estimations, and possibly unsafe seismic predictions, for the studied bridge. For the 2D model with nominal shear stiffnesses KX= KY (2D-NOM), for example, predictions of maximum reaction forces would result in discrepancies up to ≈11%, compared to the 2D-OPT model.

Tab. 3 – Seismic analysis of Dogna bridge. Comparison between maximum reaction forces FX, FY, FZ on (i) single support and (ii) single isolator, respectively. X= transversal direction; Y= longitudinal direction; Z= vertical direction. ∆= 100×(F(ii) – F(i))/F(i); ∆2D= 100×(F(OPT) – F(NOM))/F(NOM).

Traditionally supported bridge (i) Isolated bridge (ii) Isolated bridge (ii) Δ2D [%] 2D-FIX 2D-NOM 2D-OPT

[kN] [kN] [kN] [kN] Δ [%]

FX Pier �978 8�7 -87.� 841 -87.9 3.1

Abutment 31�1 844 -73.3 829 -73.8 1.8

FY Pier - 7�� - 858 - -10.7

Abutment 85�2 911 -89.4 878 -89.7 3.8

FZ Pier 15101 9980 -33.9 10304 -31.8 -3.1

Abutment 10095 9422 -�.7 9505 -�.0 -0.9

Conclusions. In the paper, dynamic characterization of the base-isolated Dogna bridge was carried out using harmonic vibration tests and FE analyses. A refined 3D FE-model of the bridge was firstly calibrated, based on natural frequencies and modal shapes extracted from

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FRFs. A theoretical procedure based on one-dimensional analytical models was also developed, to estimate the elastic stiffness of isolators. Investigations resulted in an identified elastic stiffness value about 2.2 times the expected nominal value for the used seismic devices. Based on 3D model updating, a computationally efficient 2D model of the bridge was also carried out. The optimized 2D model resulted in good agreement with experimental modal predictions and static truck-load test measurements, hence suggesting its use as baseline FE-model for further diagnostic investigations and monitoring programs. Finally, seismic analysis of 2D model under earthquake ground motions was compared to that of the same bridge with traditional restraints. Predictions clearly highlighted the marked efficiency and benefits of base isolation techniques, as well as the effects of improper FE-model calibrations for structural monitoring programs.ReferencesAlessandrini F., Fedrigo D., Coccolo A. (2009). Il nuovo ponte sismicamente isolato sul fiume Fella a Dogna:

validazione del progetto strutturale a seguito di prove dinamiche. Ingegneria Sismica XXVI(4): 41–52.Bedon C, Morassi A (2014). Dynamic testing and parameter identification of a base-isolated bridge. Engineering

Structures 60(2): 85-99.Bonessio N., Lomiento G., Benzoni G. (2012). Damage identification procedure for seismically isolated bridges,

Structural Control and Health Monitoring 19: 565–578.Naeim F., Kelly J. (1999). Design of Seismic Isolated Structures: From Theory to Practice, Wiley, New York.NTC2008. Norme tecniche per le costruzioni (in Italian), D.M. 14 January 2008.SAP2000 (1998). Integrated finite element analysis and design of structures. Berkeley, CA, USA: Computer and

Structures Inc.

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